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Contract No. DE-AC36-08GO28308

Novel Power Electronics ThreeDimensional Heat Exchanger

Preprint

K. Bennion, J. Cousineau, J. Lustbader, 

and S. Narumanchi

Presented at ITherm 2014

Orlando, Florida

May 27–30, 2014

Conference Paper

NREL/CP-5400-61041

August 2014
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Novel Power Electronics Three-Dimensional Heat Exchanger

Kevin Bennion, Justin Cousineau, Jason Lustbader, Sreekant Narumanchi

National Renewable Energy Laboratory

1617 Cole Blvd

Golden, CO, U.S.A., 80401

Email: kevin.bennion@nrel.gov

ABSTRACT

Electric-drive systems, which include electric machines 

and power electronics, are a key enabling technology to meet 

increasing automotive fuel economy standards, improve 

energy security, address environmental concerns, and support 

economic development. Enabling cost-effective electric-drive

systems requires reductions in inverter power semiconductor 

area, which increases challenges associated with heat removal.

In this paper, we demonstrate an integrated approach to the 

design of thermal management systems for power 

semiconductors that matches the passive thermal resistance of 

the packaging with the active convective cooling performance 

of the heat exchanger. The heat exchanger concept builds on 

existing semiconductor thermal management improvements 

described in literature and patents, which include improved 

bonded interface materials, direct cooling of the 

semiconductor packages, and double-sided cooling. The key 

difference in the described concept is the achievement of high 

heat transfer performance with less aggressive cooling 

techniques by optimizing the passive and active heat transfer 

paths. An extruded aluminum design was selected because of 

its lower tooling cost, higher performance, and scalability in 

comparison to cast aluminum. Results demonstrated a 102%

heat flux improvement and a package heat density 

improvement over 30%, which achieved the thermal 

performance targets.

KEY WORDS: power electronics, thermal management, 

electric drive, water-ethylene glycol, aluminum extrusions, 

electric vehicle, inverter, heat exchanger

NOMENCLATURE

A area

Al aluminum

AlN aluminum nitride

CFD computational fluid dynamics

cp coolant specific heat

Cu copper

DBC direct-bond-copper

FEA finite element analysis

HX heat exchanger

IGBT insulated gate bipolar transistor

liquid coolant mass flow

R thermal resistance

Si silicon

T temperature

WEG water-ethylene glycol

Greek symbols

ε heat exchanger effectiveness

Subscripts

b heat exchanger base

c coolant

i inlet

th,hx heat exchanger thermal resistance

INTRODUCTION

This paper describes a heat exchanger design integrated 

with the power semiconductor packages to achieve high 

thermal performance while enabling less aggressive and 

potentially lower cost convective cooling technologies 

meeting automotive requirements. The passive and active 

cooling elements of the design were refined through the use of 

software tools for thermal finite element analysis (FEA) and 

computational fluid dynamics (CFD) run on a high 

performance computer cluster to enable batch runs over many 

design variables. Multiple levels of model validation were 

performed, including a single channel analytical model to 

check CFD results, full conjugate heat transfer CFD to verify 

FEA optimization results, and experimental validation. This

work focused on application of the design to power 

semiconductor packages for automotive electric traction drive 

inverter applications.

Motivation

Electric-drive systems for vehicle propulsion enable 

technologies critical to meeting challenges for energy, 

environment, and economic security. According to the 

Transportation Energy Data Book, 70% of the total petroleum 

used in the United States is consumed by the transportation 

sector, which is dependent on petroleum as an energy source

[1]. The President's 2013 Climate Action Plan highlights that 

28% of greenhouse gas emissions were due to the 

transportation sector, and it emphasizes energy efficiency and 

technology developments to reduce emissions and increase 

energy security [2], [3]. Electric-drive systems, which include 

electric machines and power electronics, are a key enabling 

technology for advanced vehicle propulsion systems that 

reduce the dependence of the U.S. transportation sector on 

petroleum. To penetrate the market, however, these electricdrive technologies must enable vehicle solutions that are 

economically justifiable. Reducing the cost of power 

electronics is seen as a key enabler to electric propulsion [4].

To reduce power electronics cost, weight, and volume, the 

semiconductor area must be reduced. As components of the 

electric-drive system are made smaller, lighter, and more cost 

effective, heat removal becomes an increasing challenge. The 

vehicle must achieve the performance requirements as it 

operates within its thermal limitations. To successfully 

integrate advanced power electronics concepts into vehicle 

applications, the thermal limitations of the semiconductor 
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devices must be addressed. Increasing the semiconductor heat 

flux (heat per semiconductor area) reduces the required 

semiconductor area for the same power level. The reduced 

semiconductor area leads to reductions in other packaging and 

interconnect materials which further reduces cost, weight, and 

volume. This work is supported through the Department of 

Energy’s Vehicle Technologies Office in partnership with 

U.S. DRIVE to reduce cost through improved thermal 

management [5], [6].

Current Approaches to Problem

Conventional approaches to cooling power semiconductors 

are limited for vehicle applications because of the thermal 

resistance of the passive stack of the package with typical 

thermal interface material (e.g. grease) between the heat 

spreader and the heat exchanger (see Fig. 1) [7]. Methods to 

improve the passive stack thermal resistance include:

improving the thermal properties of thermal interfaces, 

eliminating thermal interfaces such as the grease layers by 

integrating the heat exchanger into the package (referred to as 

direct cooling), and employing double-sided cooling.

Fig. 1 Conventional power electronics semiconductor thermal 

stack with heat spreader and thermal interface material 

between the heat spreader and the heat exchanger surface.

Improvements to large area thermal interfaces can have a 

significant impact on improving the passive stack thermal 

resistance [7]. Research into developing and characterizing the 

performance and reliability of large area attach methods is

critical in reducing the passive stack thermal resistance [8], 

[9]. The cooling approach described in this paper benefits 

from interface material research, and future work could utilize 

improvements in large area bonding methods.

Double-sided cooling improves the power per die area of 

power semiconductors by increasing the available area for 

heat transfer. The benefits of double-sided cooling and 

example implementations are summarized in the following 

sources [10]–[18]. Double-sided cooling requires a passive 

stack cooling structure and electronics packaging capable of 

cooling both sides of the semiconductor package. This work 

attempts to develop a cooling approach that is compatible with 

single-sided and double-sided cooling. 

Another approach to improving the power capability of 

power semiconductor devices is direct cooling. Through direct 

cooling, elements within the package with higher thermal 

resistance are removed from the heat transfer path by 

removing layers and applying the convective cooling closer to 

the heat source. The benefits of direct cooling as highlighted 

by Skuriat and Johnson [19] include reduction of package 

thermal resistance through removal of material layers such as 

grease, and conventional cold plates. Also, Skuriat and 

Johnson mention fewer interfaces with different thermal 

expansion coefficients lead to potential improvements in 

thermal reliability.

Multiple methods have been used or proposed for direct 

cooling of the semiconductor. One approach involves 

removing the grease thermal interface material and directly 

cooling the base plate or heat spreader of the power 

semiconductor module. Examples include pins molded or 

machined into the heat spreader [20]–[22], folded fins bonded 

to the heat spreader [23], or jet impingement onto the heat 

spreader [24]–[26]. Beside cooling the base plate, other 

examples include directly cooling the substrate material or the 

metalization layers on the substrate, such as direct bond 

copper (DBC), using jets [27]–[30], or through the use of 

microchannels [31], [32]. Also, microchannels within the 

silicon and silicon carbide layers have been investigated [33]–

[35]

As layers are removed within the package to reduce the 

passive stack thermal resistance, typically the available 

footprint area to which cooling can be applied is also reduced.

The reduction in area necessitates higher performing or more 

aggressive cooling technologies. For this reason, the cooling 

performance must be matched to the package design to get the 

optimal integrated solution [10], [36]. The more aggressive 

cooling can potentially be difficult to manufacture in highvolume applications required for automotive systems, which 

shifts more cost into the thermal management system. The 

increased cost can reduce or cancel the cost savings of reduced 

silicon area.

When developing a package that is directly cooled, other 

thermal bottlenecks could be introduced in the power 

electronics system that might reduce the effectiveness of the 

proposed thermal design. Possible examples include the 

following items. The thermal performance of the power 

semiconductor could degrade by increasing the thickness of 

low thermal conductivity materials. An integrated package 

requires methods of bonding or attaching the cooling interface 

to the package that are capable of withstanding the potential 

thermal stresses. The bonded interface can lead to thermal 

bottlenecks. Unlike applications that have steady continuous 

loads, electric-drive systems in vehicle applications experience 

significant dynamic changes in loading. The effective thermal 

capacitance of the packaging structure provides some support 

in reducing the magnitude of temperature cycles [37], [38] and 

the impact on reliability. Also, improvements in 

semiconductor thermal management often focus on the 

semiconductor and neglect other temperature sensitive power 

electronics components. For example, the bus bars can 

conduct heat to other components [39]. Finally, another 

concern is related to sealing of the cooling system to prevent 

coolant leaks into the electronics [21].

APPROACH

We optimized the passive and active thermal design 

components for an inverter application. The passive thermal 

design focused on the mechanical structure of the 

semiconductor package and heat exchanger. The thermal 

performance of the passive components is a function of the 

layer material properties, interfaces, and the layer material 

geometries. The passive thermal performance is also affected 
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by the technology selected for the active thermal design. The 

active thermal design focused on the convective cooling 

interface between the heat exchanger and the cooling fluid. A 

cost effective solution requires matching the passive thermal 

design with the active cooling design. The approach to

designing an integrated passive and active thermal 

management system for power semiconductor modules 

included developing design specifications, designing the 

prototype, and testing the concept experimentally.

Design Specifications

The design specifications emphasize thermal management

because the project focused on thermal management of the 

power semiconductor devices and not development of a full 

inverter. The achievable power per silicon area directly 

impacts the cost of power electronics, which is limited by the 

ability to remove heat. For this reason, the heat removed from 

the Insulated Gate Bipolar Transistor (IGBT) per footprint 

area of the IGBT was the primary metric to illustrate the 

potential relative cost benefit. The design target was to double 

the IGBT heat flux as compared to a best-in-class commercial 

module.

A cost effective thermal management approach also 

required compatibility with established high volume 

manufacturing methods used within automotive applications.

Examples include aluminum castings, brazed fins, and 

aluminum extrusions. The selected design focused on using 

aluminum extrusion techniques to reduce fabrication cost 

through adherence to industry supplied guidelines and 

reducing voids or hollow channels within the part. The design 

also emphasized a modular, scalable, and flexible design. A 

modular design allows the package to be reused and has 

broader applications. For example, the goal was to develop a 

thermal design that could scale up in power requirements 

without the need for a complete redesign of a heat exchanger 

casting. Another reason an aluminum extrusion process was 

selected is because aluminum extrusions can have higher 

thermal conductivity and lower cost as compared to cast 

structures if designed properly for the targeted application.

In addition to cost, the volume of the power electronics is 

another critical design goal because of space limitations within 

a vehicle. The volume of the power electronics not only 

depends on the size of the components but also on how 

densely the components can be packaged while still removing 

heat. The volume specifications were described in terms of the 

power module heat density. The power module heat density is 

defined as the total heat removed divided by the total volume 

of the semiconductor power module packaging and heat 

exchangers. The design target was to maintain or exceed

power module heat density as compared to a best-in-class 

commercial module.

Additional design constraints focused on factors to reduce 

failure modes associated with the thermal management 

system. The minimum coolant channel dimensions must allow 

a 1 mm particle to pass through the cooling system without 

clogging. Eliminating internal fluid seals around the power 

semiconductor devices reduces the impact of fluid leaks 

within the electronics. Maintaining or increasing thermal 

capacitance of the module reduces the sensitivity to 

temperature cycling. Finally, the design should also be 

compatible with cooling or isolating other temperature 

sensitive components within the power electronics. For 

example, bus bar cooling could thermally isolate temperature 

sensitive components and suppress heat transfer from the 

motor to the inverter [39].

To consistently evaluate progress towards the design 

specifications, the following analysis assumptions were 

applied. The IGBT and diode operating losses or heat loads

were proportioned with an IGBT-to-diode heating ratio of 

three to one. The maximum semiconductor (IGBT, diode) 

temperature was limited to 150°C to calculate the maximum 

heat flux removal rates. Liquid cooling with a 50/50 by mass 

mixture of water-ethylene glycol (WEG) was assumed. The 

design flow rate at the inverter system level was 1.67e-4 m3

/s 

(10 L/min) with a 70°C inlet coolant temperature.

The package structure was limited to conventional in-use 

materials as shown in Fig. 2. However, removal of layers was 

allowed. The design was also required to be compatible with 

single-sided and double-sided cooling. Focusing on existing 

packaging methods and materials prevented the introduction 

of unknown costs through new materials or manufacturing 

processes. The material properties are listed in Table 1.

Fig. 2 Schematic of package layers. 

Table 1. Material properties included in analysis

Layer Material Thermal 

Conductivity 

[W/m-K]

A IGBT/Diode Si 92

B Bonded Interface Solder 51

C Metalization Cu 394

D Substrate AlN 140

E Metalization Cu 394

F Bonded Interface Solder 51

G Heat Spreader Cu or

Aluminum (Al) 6063

394 or

210

H Thermal Interface Grease 1.5

I Heat Sink Al 6061 154

Design Approach

We identified the appropriate passive and active thermal 

designs to meet performance targets through comparison to 

selected baseline packages. We optimized the selected passive 

thermal design and identified the active thermal design targets.

The primary commercial system benchmark was the power 

module used within the LS 600h hybrid system, which is 

described in detail in [11], and the thermal performance is 

based on analysis described in [10]. The LS 600h package was 

selected because of its best-in-class performance as 

benchmarked through the Vehicle Technology Office 

Advanced Power Electronics and Electric Motors activity. The 

design is also similar to recent hybrid systems for lower cost 

vehicles. Potential designs were also compared against direct 
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cooling approaches for a direct-cooled base plate and directcooled DBC. The direct-cooled DBC performs best at low 

thermal resistance as seen in Fig. 3.

The targeted minimum heat exchanger cooling resistance 

was between 100 to 200 mm2

-K/W, shown in Fig 3. This 

range represents typical channel flow cooling with WEG. The 

ability to enable higher resistance, leads to potential cost 

reductions in the heat exchanger fabrication and increased 

system robustness. Fig. 3 highlights that the existing direct 

cooling approaches do not achieve the desired IGBT heat flux 

improvement within the targeted active cooling performance 

region. The targeted active cooling performance region in Fig.

3 is defined as the heat exchanger thermal resistance defined 

below in Equation 1. The heat exchanger thermal resistance

(Rth,hx) is based on the cooled base area (Ab), heat exchanger 

effectiveness (ε), coolant mass flow ( ), and coolant specific 

heat (cp) [40].

(1)

Fig. 3 Package comparison against targets assuming ability for 

double-sided cooling on all configurations.

Fig. 4 Conceptual drawings of integrated heat spreader and 

heat exchanger system to enhance heat spreading and 

cooling performance.

After establishing the active cooling target, the task of 

designing a passive thermal system was initiated to meet the 

performance targets. The heat spreader geometry was 

optimized in FEA for the targeted active cooling convection 

coefficient range. The final passive cooling design is 

illustrated in Fig. 4. The design removes the heat sink and

grease thermal interface layers (layers “H” and “I” in Fig. 2)

to directly cool the heat spreader. The expanded heat spreader 

increases the cooled area by cooling both the back side of the 

heat spreader and the exposed package side of the heat 

spreader. Unlike conventional cooling approaches that are 

dominated by one dimensional heat transfer to a heat 

exchanger, the design enables heat to spread in all directions 

to multiple cooling zones within the heat spreader. For this 

reason it is described as a three-dimensional heat exchanger.

The design process required an iterative procedure to meet the 

IGBT heat flux and module heat density targets within the 

minimum active cooling performance target. For each design, 

the volumetric IGBT and diode heat loads were adjusted until 

the maximum IGBT temperature limit of 150°C was reached. 

Efforts were made to maintain simple geometries to enable the 

use of aluminum extrusions based on industry input, and the 

general concept was described in a patent [41].

Once the passive thermal stack was designed, the next step 

involved designing the active cooling surface through both 

FEA and CFD analysis. CFD modeling tools enabled the 

evaluation of multiple cooling channel geometries.

Throughout the analysis the models were checked for mesh 

independence and conservation of mass and energy. In the 

absence of preexisting experimental data for the preliminary 

design, other methods of cross checking the model accuracy 

were employed. The channel convection coefficients were 

compared against analytical estimates, while model 

decomposition of full conjugate heat transfer CFD simulations

provided preliminary tests for modeling errors. The full 

conjugate heat transfer CFD conduction and convection results 

were checked independently. Separating the conduction and 

convection results provided comparisons to simplified models 

with boundary conditions calculated from the full conjugate 

heat transfer CFD simulations. Mapping the convection 

coefficient obtained from the CFD simulation to an FEA 

model enabled comparison of the solid-only heat conduction 

model accuracy of the CFD simulation.

The final design was developed with an emphasis on ease

of manufacture based on industry input to enable less 

aggressive cooling while achieving the thermal performance 

targets. A few of the key design features relevant to the 

aluminum extrusion process are highlighted in Fig. 5. The 

coolant channels and fins were made as large as possible to 

reduce channel clogging concerns and to increase ease of 

extrusion. The open “E-fin” shape within the connected 

channels significantly reduces the cost of the extrusion by 

reducing the number of internal voids and the extrusion 

complexity. The rounded corners and minimum wall thickness 

follow recommended design guidelines for aluminum 

extrusions based on input from Sapa Extrusions North 

America. The final prototype design is shown in Fig. 6 with 

flanges on the inlet and outlet for attachment to the 

experimental test setup. A heater is also shown in the figure 

mounted to the heat exchanger for thermal testing.

Fig. 5 Sample parameters from design guidelines to reduce 

fabrication cost.
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Fig. 6 Prototype design with flanges for connection to 

experimental test setup and a heater for thermal testing.

Experimental Approach

Once the design was selected, three heat exchangers were 

built. Two heat exchangers representing the design were built 

with Al alloy 6061 and Al 6063. The third heat exchanger,

representing a baseline with a single cooling zone beneath the 

package, was made with Al 6061. The baseline and design 

heat exchanger geometries are shown in Fig. 7. The intended 

target material was Al 6063 because it is commonly used in 

extruded shapes, and it has a higher thermal conductivity 

relative to Al 6061. Al 6061 was used initially to build two 

heat exchangers for comparing the design against the baseline 

geometry and validate CFD modeling. Al 6061 was selected

for the initial tests because of availability of the bulk material.

Later a third prototype was manufactured using Al 6063 

provided by Sapa Extrusions North America, which was used 

to confirm the expected performance benefit over Al 6061.

The fabricated heat exchangers and materials are summarized 

in Table 2.

Fig. 7 Fabricated heat exchanger geometries showing baseline 

(top) and design (bottom).

Table 2. Tested heat exchangers and materials

Geometry Material

Baseline Al 6061

Design Al 6061

Design Al 6063

To test the prototype heat exchangers and validate the 

modeling results, we designed and built the experimental 

apparatus highlighted in Fig. 8. The prototype heat exchanger 

was installed between two manifolds of similar design. The 

coolant flow develops in the manifold prior to entering the 

heat exchanger channels to isolate potential flow variations 

caused by the inlet and outlet connections to the fluid test 

bench. The manifolds also provide locations for measuring 

inlet and outlet temperatures and pressures. The manifolds 

were designed using CFD models to determine the

temperature and pressure measurement locations. The pressure 

taps are located upstream of the temperature measurements.

Fig. 8 Heat exchanger installed in test fixture with inlet and 

outlet temperature and pressure measurements.

As shown in Fig. 6, a ceramic heater was used to simulate 

a power semiconductor device heat load. The heater was 

mounted to a copper spreader with epoxy (Fig. 9). The copper 

spreader was bonded to the heat exchanger with a

thermoplastic material made by Btechcorp (referred to in this 

paper as Btech). The copper spreader was instrumented with 

thermocouples in two locations (Fig. 10). The thermocouples 

were soldered into machined grooves in the copper spreader to 

ensure good thermal contact. To prevent damage to the 

heating element, the heater also included a temperature 

measurement.

The measurement uncertainties of temperature and 

pressure were estimated according to the 95% uncertainty 

model (U95) [42], including both systematic and random 

errors. The type K thermocouples were calibrated over the 

operating temperature range using a reference (NIST 

traceable) probe and calibration bath. The total U95 confidence 

interval for the temperature measurements were within 

±0.13°C (~0.4% of average measurement). The inlet and 

outlet pressures were monitored with a U-tube manometer 

with a U95 of ±14 Pa (~1% of average measurement).

Fig. 9 Heater assembly layers and attachment to heat 

exchanger.

Fig. 10 Thermocouple locations in the copper spreader.
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Each of the heat exchangers listed in Table 2 was tested 

experimentally to verify the simulated thermal and fluid 

characteristics of the heat exchanger from the CFD simulation 

results. The model validation experiments were performed 

with a WEG mixture of 50%/50% by volume. The heat 

exchangers were tested at 0.033 kg/s, 0.067 kg/s, and 0.1 kg/s

(1.9 L/min, 3.8 L/min, and 5.7 L/min). The tests for each heat 

exchanger included six repetitions for each test condition. Due 

to equipment limitations and safety, the inlet coolant 

temperature was set to 30°C for each test. The Specific gravity 

of the WEG was checked prior to starting tests to ensure the 

fluid properties remained consistent and similar to the 

simulated coolant properties.

The epoxy layer and the Btech layer illustrated in Fig. 11

represented unknown thermal resistance within the passive

thermal stack. The values for these thermal resistances were 

estimated by matching the temperature measurements at the 

lowest flow rate to the model results at the same fluid flow 

rate and fluid properties. The process for determining the 

effective thermal conductivity of the epoxy and Btech layers 

followed the steps listed below.

1. The convection coefficients were mapped from a single 

flow rate CFD model to the channels in the FEA model.

2. The effective thermal conductivity of the epoxy and 

Btech layers were solved iteratively for a single test by 

matching the temperature change between the heater to 

copper spreader and the temperature change between 

the copper spreader to the inlet coolant temperature

(Fig. 11).

3. The effective thermal resistance values for the epoxy 

and Btech layers were applied to the CFD model and 

validated against experimental results for other flow 

rates.

Fig. 11 Interface thermal resistance estimation for epoxy and 

Btech bonding layers.

Using the above approach for each of the prototype heat 

exchangers, we validated the CFD model results against the 

experimental data. The corner thermocouple temperature 

measurement in the copper spreader was compared against a 

corresponding point within the CFD model. Fig. 12 compares 

the difference between the experimental temperature and CFD 

model temperature within the copper spreader block. The 

magnitude of the difference between the model and the 

experimental results was found to be within about 2% over all 

of the flow rates and heat exchangers tested. This accuracy 

was within the targeted magnitude limit of 10%.

Fig. 13 compares the experimentally measured pressure 

drop with the CFD model results. The magnitude of the 

difference between the model and the experimental results was 

within about 9% across all flow rates and heat exchangers that 

were tested. With the analysis, it was found that the pressure 

drop was affected by small variations in the channel geometry. 

We used a high-resolution digital microscope to measure the 

manufactured channel dimensions for each heat exchanger and 

applied the averaged results to the CFD model geometries 

corresponding to each heat exchanger. This substantially

improved the agreement of the CFD simulation results with 

the experimental results for all flow rates. The difference 

between the model and experimental results was within the 

targeted maximum magnitude of 10%.

Fig. 12 Temperature validation for all tested heat exchangers 

across all flow rates.

Fig. 13 Pressure validation for all tested heat exchangers 

across all flow rates.

RESULTS

The design met or exceeded the thermal system level (heat 

exchanger and semiconductor package) performance targets

within the established design constraints and evaluation 

criteria. The combined semiconductor package model and 

validated heat exchanger model confirmed the expected 

system level performance impacts. The system level impacts 

of the heat exchanger design were determined through 

simulation using the validated heat exchanger CFD model. 

The system performance as compared to the baseline reference 

package is shown in Fig. 14. The primary goal of the research 

was to enable high thermal performance (IGBT heat flux) with 

less aggressive cooling (higher heat exchanger side thermal 
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resistance). Early in the design process, a targeted cooling 

performance range was identified as highlighted in Figs. 3 and

14, which was selected from past experience related to power 

electronics cooling with WEG. As seen in Fig. 14, the final 

design achieves the targeted improvement of a 100% 

improvement in IGBT heat flux within the targeted cooling 

performance region.

Fig. 14 Thermal FEA comparing IGBT heat flux of heat 

spreader design showing the original target cooling 

performance range and the final design.

Fig. 15 compares the design against the LS 600h reference 

package and the baseline aluminum heat exchanger described 

in Fig. 7. The comparison in Fig. 15 is at the equivalent flow 

rate for the LS 600h package (0.0086 kg/s per side of 

package), which is based on referenced modeling and analysis 

efforts [10]. At an equivalent flow rate, the design improves 

the IGBT heat flux by a factor of 1.6 with a slight 

improvement in package heat density. When the flow rate is 

held constant, the pressure drop across the design heat 

exchanger is significantly lower than the LS 600h reference. 

With the improved heat transfer and reduced pressure drop, 

the coefficient of performance for the design improves by a 

factor of 7.9 as compared to the reference LS 600h package.

Fig. 15 Model results for design and baseline comparison at 

LS 600h package flow rate.

When evaluated at the target system flow rate of 1.67e-4 

m3

/s (10 L/min) in the parallel configuration shown in Fig. 16, 

the design meets the performance targets defined at the 

beginning of the work. Fig. 16 shows the results with 1/6 of 

the system coolant flowing into each package. The IGBT heat 

flux improved by 102% and the package heat density 

improved by 32%. The design heat exchanger also exceeds the 

performance of the baseline at the same fluid parasitic power.

Fig. 16 Model results for design at target flow and baseline at 

equivalent parasitic fluid power.

As mentioned previously, the design was based on a 

parallel arrangement of the heat exchangers as shown in Fig.

16 with one package or module per cooling branch. 

Preliminary investigations were also performed to investigate 

the impact of series connections on the effective convection 

coefficient within the channels. Fig. 17 compares the average 

convection coefficient within the channels for three separate 

heat exchangers connected in series. As seen in Fig. 17, the 

convection coefficient drops as the flow passes through the 

first heat exchanger. The convection coefficient appears to 

stabilize as the flow passes through the second and third heat 

exchangers. The drop in the convection coefficient would need 

to be investigated as part of future work depending on the 

intended application. We attributed the drop in convection 

coefficient to the flow becoming fully developed by the time it 

reaches the second heat exchanger.

Fig. 17 Effective channel convection coefficient in each 

section.

SUMMARY AND CONCLUSIONS

This work focused on the development and demonstration 

of a concept to enable less aggressive cooling while improving 

the system-level semiconductor package thermal performance

to improve the capability of power semiconductor devices. 
8

This report is available at no cost from the National Renewable Energy Laboratory (NREL) at www.nrel.gov/publications.

The improved power capability directly relates to improving 

the cost per power of the power electronics. The performance 

of the design was compared against two baseline packages of 

power semiconductor devices. The design doubled the IGBT 

heat flux capability relative to the LS 600h reference with 

significant improvements in the coefficient of performance. 

The LS 600h package was selected as the reference because of 

its best-in-class performance of the modules benchmarked 

through the Vehicle Technology Office’s Advanced Power 

Electronics and Electric Motors activity. The design also 

outperformed the baseline direct-cooled base plate 

configuration with similar extruded fin geometry.

The design approach built upon a recently issued patent

[41] to enable high performance with less aggressive cooling. 

Future work could extend the concept to enable other cooling 

technologies with less effective coolants such as air or 

automatic transmission fluid. Directly cooling the power 

electronics with air or automatic transmission fluid could 

eliminate the dedicated WEG loop for cooling the power 

electronics and lead to potential cost reductions and 

elimination of cooling system components. While the design

included efforts to enable cooling of bus bars or other 

temperature sensitive components, additional design analysis 

and refinements are needed in these areas specific to target 

applications. Also, research into improved large area bonding 

materials with good thermal performance and temperature 

cycling ability is necessary. The current work assumed a 

bonded interface between the DBC and heat spreader with 

equivalent thermal performance as solder. Additional work is 

needed to evaluate the reliability of high thermal performance 

bonded interfaces between the DBC and aluminum heat 

spreader used in this proposed design.

This work developed an integrated design process and 

applied the process to design an integrated power module heat 

exchanger, which builds on existing developments in bonded 

interfaces, direct cooling, and double sided cooling. The 

specific design can go through additional refinement for 

specific power electronics applications and this work will be 

included in future industry collaborations related to cooling 

power electronics.

ACKNOWLEDGMENTS

The authors would like to acknowledge the support 

provided by Susan Rogers and Steven Boyd, Technology 

Development Managers for Advanced Power Electronics and 

Electric Motors, Vehicle Technologies Office, U.S. 

Department of Energy Office of Energy Efficiency and 

Renewable Energy. The significant contributions of Sapa 

Extrusions North America to the project are acknowledged.

REFERENCES

[1] S. Davis, S. Diegel, and R. Boundy, Transportation 

Energy Data Book, 32nd ed. Oak Ridge National 

Laboratory, 2013.

[2] “The President’s Climate Action Plan,” Jun-2013. 

[Online]. Available: 

http://www.whitehouse.gov/sites/default/files/image/pre

sident27sclimateactionplan.pdf. [Accessed: 21-Aug2013].

[3] The White House, “President Obama’s Plan to Fight 

Climate Change,” 25-Jun-2013. [Online]. Available: 

http://www.whitehouse.gov/share/climate-action-plan. 

[Accessed: 21-Aug-2013].

[4] “EV Everywhere Grand Challenge Blueprint.” [Online]. 

Available: 

http://www1.eere.energy.gov/vehiclesandfuels/electric_

vehicles/pdfs/eveverywhere_blueprint.pdf. [Accessed: 

08-Aug-2013].

[5] “Vehicle Technologies Office: Advanced Power 

Electronics and Electrical Machines.” [Online]. 

Available: 

http://www1.eere.energy.gov/vehiclesandfuels/technolo

gies/electronics/index.html. [Accessed: 10-Sep-2013].

[6] U.S. Drive, “Electrical and Electronics Technical Team 

Roadmap,” Partnership Plan, Roadmaps, and Other 

Documents, Jun-2013. [Online]. Available: 

http://www1.eere.energy.gov/vehiclesandfuels/pdfs/pro

gram/eett_roadmap_june2013.pdf. [Accessed: 10-Sep2013].

[7] S. Narumanchi, M. Mihalic, K. Kelly, and G. Eesley, 

“Thermal Interface Materials for Power Electronics 

Applications,” in 11th Intersociety Conference on 

Thermal and Thermomechanical Phenomena in 

Electronic Systems (ITherm), Orlando, FL, 2008, pp. 

395–404.

[8] D. DeVoto, P. Paret, S. Narumanchi, and M. Mihalic, 

“Reliability of Bonded Interfaces for Automotive Power 

Electronics,” in ASME 2013 International Technical 

Conference and Exhibition on Packaging and 

Integration of Electronic and Photonic Microsystems, 

Burlingame, CA, 2013.

[9] T. G. Lei, J. N. Calata, G.-Q. Lu, X. Chen, and S. Luo, 

“Low-Temperature Sintering of Nanoscale Silver Paste 

for Attaching Large-Area Chips,” IEEE Trans. Compon. 

Packag. Technol., vol. 33, no. 1, pp. 98–104, 2010.

[10] K. Bennion and K. Kelly, “Rapid Modeling of Power 

Electronics Thermal Management Technologies,” in 

IEEE Vehicle Power and Propulsion Conference 

(VPPC), Dearborn, MI, 2009, pp. 622–629.

[11] T. Burress, C. Coomer, S. Campbell, A. Wereszczak, J. 

Cunningham, L. Marlino, L. Seiber, and H.-T. Lin, 

“Evaluation of the 2008 Lexus LS 600H Hybrid 

Synergy Drive System.” Oak Ridge National 

Laboratory. ORNL/TM-2008/185, Jan-2009.

[12] H. Yasui, H. Ishiyama, M. Inagaki, K. Mamitsu, and T. 

Kikuchi, “Power Control Unit for High Power Hybrid 

System,” presented at the 23rd International Electric 

Vehicle Symposium, Anaheim, California, 2007.

[13] Y. Sakai, H. Ishiyama, and T. Kikuchi, “Power Control 

Unit for High Power Hybrid System,” SAE 

International, Warrendale, PA, 2007-01-0271, Apr. 

2007.

[14] M. Inagaki and M. Shirai, “Stacked Type Cooler,” U.S. 

Patent 7571759,11-Aug-2009.

[15] D. Harada and H. Ishiyama, “Power Converter and 

Semiconductor Device Mounting Structure,” U.S. 

Patent Application 2006/0284308 A1.
9

This report is available at no cost from the National Renewable Energy Laboratory (NREL) at www.nrel.gov/publications.

[16] R. S. Taylor, M. A. Jeter, E. W. Gerbsch, and J. J. 

Ronning, “Electronics Assembly Having Multiple Side 

Cooling and Method,” U.S. Patent 7295433,13-Nov2007.

[17] M. Inagaki, “Easily Assembled Cooler,” U.S. Patent 

8151868,10-Apr-2012.

[18] M. O’Keefe and K. Bennion, “A Comparison of Hybrid 

Electric Vehicle Power Electronics Cooling Options,” in 

IEEE Vehicle Power and Propulsion Conference 

(VPPC), Arlington, TX, 2007, pp. 116–123.

[19] R. Skuriat and C. M. Johnson, “Direct Substrate 

Cooling of Power Electronics,” in 2008 5th 

International Conference on Integrated Power Systems 

(CIPS), Nuremberg, Germany, 2008, pp. 1–5.

[20] F. Nishikimi and K. Nakatsu, “Power Inverter,” U.S. 

Patent Application 2012/0170217 A1.

[21] T. Kurosu, K. Sasaki, A. Nishihara, and K. Horiuchi, 

“Packaging Technologies of Direct-Cooled Power 

Module,” in International Power Electronics 

Conference (IPEC), Sapporo, Japan, 2010, pp. 2115–

2119.

[22] R. Bayerer, “Advanced Packaging Yields Higher 

Performance and Reliability in Power Electronics,” 

Microelectron. Reliab., vol. 50, pp. 1715–1719, 2010.

[23] T. Matsuo, H. Hamada, A. Nishihara, and M. Musou, 

“Power Inverter,” U.S. Patent 8072760,06-Dec-2011.

[24] E. M. Dede, “Jet Impingement Heat Exchanger 

Apparatuses and Power Electronics Modules,” U.S. 

Patent 8199505,12-Jun-2012.

[25] E. M. Dede, “Experimental Investigation of the Thermal 

Performance of a Manifold Hierarchical Microchannel 

Cold Plate,” in ASME 2011 Pacific Rim Technical 

Conference and Exhibition on Packaging and 

Integration of Electronic and Photonic Systems 

(InterPACK), Portland, OR, 2011, vol. 2, pp. 59–67.

[26] S. Narumanchi, M. Mihalic, G. Moreno, and K. 

Bennion, “Design of Light-Weight, Single-Phase 

Liquid-Cooled Heat Exchanger for Automotive Power 

Electronics,” in 13th IEEE Intersociety Conference on 

Thermal and Thermomechanical Phenomena in 

Electronic Systems (ITherm), San Diego, CA, 2012, pp. 

693 –699.

[27] G. Moreno, S. Narumanchi, T. Venson, and K. Bennion, 

“Microstructured Surfaces for Single-Phase Jet 

Impingement Heat Transfer Enhancement,” J. Therm. 

Sci. Eng. Appl., vol. 5, no. 3, Jun. 2013.

[28] V. Hassani, A. Vlahinos, and D. Bharathan, “Low 

Thermal Resistance Power Module Assembly,” U.S. 

Patent 7190581,13-Mar-2007.

[29] K. Le, T. G. Ward, B. S. Mann, E. P. Yankoski, and G. 

S. Smith, “Power Electronics Substrate for Direct 

Substrate Cooling,” U.S. Patent 8169779,01-May-2012.

[30] E. M. Dede, J. Lee, Y. Liu, B. Robert, and S. H. Yonak, 

“Computational Methods for the Optimization and 

Design of Electromechanical Vehicle Systems,” Int. J. 

Veh. Des., vol. 58, no. 2–4, pp. 159–180, 2012.

[31] J. Schulz-Harder, “Efficient Cooling of Power 

Electronics,” in 3rd International Conference on Power 

Electronics Systems and Applications (PESA), Hong 

Kong, China, 2009, pp. 1–4.

[32] D. J. Sharar, N. R. Jankowski, and B. Morgan, 

“Thermal Performance of a Direct-Bond-Copper 

Aluminum Nitride Manifold-Microchannel Cooler,” in 

26th Annual IEEE Semiconductor Thermal 

Measurement and Management Symposium (SEMITHERM), Santa Clara, CA, 2010, pp. 68–73.

[33] C. Bower, A. Ortega, P. Skandakumaran, R. 

Vaidyanathan, and T. Phillips, “Heat Transfer in WaterCooled Silicon Carbide Milli-Channel Heat Sinks for 

High Power Electronic Applications,” J. Heat Transf., 

vol. 127, no. 1, pp. 59–65, Jan. 2005.

[34] J. A. Carter, L. A. Forster, and M. D. Stitt, “Fabrication 

and Performance of Tree-Branch Microchannels in 

Silicon Carbide for Direct Cooling of High-Power 

Electronics Applications,” in 25th Annual IEEE 

Semiconductor Thermal Measurement and Management 

Symposium (SEMI-THERM), San Jose, CA, 2009, pp. 

128–133.

[35] D. B. Tuckerman and R. F. W. Pease, “HighPerformance Heat Sinking for VLSI,” IEEE Electron 

Device Lett., vol. 2, no. 5, pp. 126– 129, May 1981.

[36] K. Bennion and G. Moreno, “Thermal Management of 

Power Semiconductor Packages — Matching Cooling 

Technologies with Packaging Technologies,” presented 

at the IMAPS 2nd Advanced Technology Workshop on 

Automotive Microelectronics and Packaging, Dearborn, 

MI, 2010.

[37] H. Ohtsuka and F. Anraku, “Development of Inverter 

for 2006 Model Year Civic Hybrid,” in Power 

Conversion Conference (PCC), Nagoya, Japan, 2007, 

pp. 1596–1600.

[38] F. Anraku, H. Ohtsuka, and T. Masuda, “Compact and 

High-Power Inverter for 2006 Model Year Civic 

Hybrid,” in The 22nd International Battery, Hybrid and 

Fuel Cell Electric Vehicle Symposium & Exposition, 

Yokohama, Japan, 2006.

[39] H. Shimizu, T. Okubo, I. Hirano, S. Ishikawa, and M. 

Abe, “Development of an Integrated Electrified 

Powertrain for a Newly Developed Electric Vehicle,” 

SAE International, Warrendale, PA, SAE Technical 

Paper 2013-01-1759, Apr. 2013.

[40] R. J. Moffat, “Modeling Air-Cooled Heat Sinks as Heat 

Exchangers,” in Twenty Third Annual IEEE 

Semiconductor Thermal Measurement and Management 

Symposium (SEMI-THERM), San Jose, CA, 2007, pp. 

200–207.

[41] K. Bennion and J. Lustbader, “Integrated ThreeDimensional Module Heat Exchanger for Power 

Electronics Cooling,” U.S. Patent 8541875,24-Sep2013.

[42] R. H. Dieck, Measurement Uncertainty: Methods and 

Applications. ISA, 2007.